Thermal Spraying Nozzle Device and Thermal Spraying System

ABSTRACT

[Subject]A thermal spraying nozzle device and a thermal spraying system are to be provided which can supply a thermal spraying material constantly and can control the state of a film or deposit. [Solution] 
     In a thermal spraying nozzle device wherein carrier gas is introduced into an inlet side of a nozzle to form a supersonic gas flow and a thermal spraying material is atomized and ejected by the gas flow, a storage section ( 4 ) for the storage of molten metal as the thermal spraying material is connected an end portion on the inlet side of the nozzle ( 2 ) through a connecting pipe, the nozzle has a throat portion ( 2   a ) for accelerating the introduced carrier gas to supersonic velocity and a divergent region ( 2   b ) formed downstream of the throat portion toward an outlet, the metal particles atomized by the supersonic gas being cooled to a solidified or semi-solidified state in the divergent region.

FIELD OF ART

The present invention relates to a thermal spraying nozzle device, aswell as a thermal spraying system, wherein with use of gas a thermalspraying material is atomized and brought into collision with a basematerial to form a film or a deposition layer.

BACKGROUND ART

Heretofore, a thermal spraying process has been known as a technique ofheating a coating material and the resulting melted or half-melted fineparticles are brought into collision at high speed with the surface of abase material to form a film.

According to this thermal spraying process, since the base material andthe film are joined together physically, a film can be formed on anymaterial insofar as the material can be melted. The film formed cansatisfy various conditions required in surface treatment, includingabrasion resistance, corrosion resistance and heat insulating property,and is therefore utilized widely in various fields.

According to a cold spraying method, a spraying material is brought intocollision with a base material as a supersonic flow together with inertgas and as it is in a solid phase without being melted or gasified, toform a film. Thus, unlike other spraying methods, the cold sprayingmethod is advantageous in that there occurs no thermal change incharacteristics of the material and that it is possible to suppressoxidation in the resulting film.

FIG. 32 shows a schematic construction of a cold spraying system.

In the same figure, high pressure gas supplied from a gas source 30 isbranched into two pipes 31 and 32. The gas as a main flow in the pipe 31is heated by a gas heater 33, while the remaining gas flowing in thepipe 32 is introduced into a powder supply unit 34.

The gas heated by the gas heater 33 is introduced into a chamber 36through a pipe 35, while the powder supply unit 34 supplies powderparticles to the chamber 36 through a pipe 37.

The gas and the powder particles are mixed together within the chamber36 and the resulting mixture passes through a convergent portion 38 aand a diffusion portion 38 b in a supersonic nozzle 38, whereby themixture is accelerated and collides as a supersonic jet flow onto a basematerial 39 (see, for example, Japanese Patent Laid-Open No.2004-76157,Patent Literature 1).

On the other hand, there also has been proposed a method wherein moltenmetal is used as a thermal spraying material and is allowed to flow inthe state of a thin film from a chamber having a slit-like outlet, thenis atomized and sprayed by a supersonic gas flow which passes in thestate of a laminar flow through a nozzle, the nozzle having a slit-likeorifice formed in the vicinity of an outlet of the nozzle (see, forexample, Japanese Patent Laid-Open No.2002-508441, Patent Literature 2).

Patent Literature 1:

-   -   Japanese Patent Laid-Open No. 2004-76157

Patent Literature 2:

-   -   Japanese Patent Laid-Open No. 2002-508441

DISCLOSURE OF THE INVENTION

However, in the former cold spraying system referred to above, sincepowder particles of a normal temperature are brought into collision witha base material and are heated locally to a temperature of not lowerthan the melting point thereof with heat resulting from plasticdeformation and thereby adhered onto the base material, a gas pressureof 1.0 to 3.0 MPa is needed for attaining a particles velocity of, say,600 m/s or more. Besides, it is necessary to pre-heat the gas up to 600°C. Thus, there is the problem that handling is difficult. Moreover, itis not easy to supply the powder particles at a constant rate.

In the latter thermal spraying system referred to above, atomization isperformed at a supersonic velocity, the nozzle is not designed foracceleration of particles and therefore it is impossible to obtain suchhigh density film or deposit as permits omission of HIP (Hot IsostaticPressing).

The present invention has been accomplished in view of theabove-mentioned problems involved in the conventional spraying systemsand provides a thermal spraying nozzle device and a thermal sprayingsystem both able to supply a thermal spraying material at a constantrate and control the state of a film or a deposit.

The thermal spraying nozzle device according to the present inventionis, in the gist thereof, a thermal spraying nozzle device whereincarrier gas is introduced from an inlet side of a nozzle to form asupersonic gas flow and a thermal spraying material is atomized andejected by the gas flow, the thermal spraying nozzle device comprising,a storage section storing molten metal as the thermal spraying materialconnected to an end on the inlet side of the nozzle through a connectingpipe, and, the nozzle having a throat portion and a divergent region ina downstream of the throat portion toward an outlet side to form thesupersonic gas flow, wherein the thermal spraying nozzle device isconfigured such that metal particles atomized by the supersonic gas floware cooled to a solidified or semi-solidified state in the divergentregion and then ejecting in a predetermined direction from the outletside of the nozzle.

Preferably, within the connecting pipe in the above thermal sprayingnozzle device, a molten metal outlet pipe is extended from the storagesection toward the center in the throat portion or the center on thedownstream side of the throat portion and an outside portion of themolten metal outlet pipe constitutes a channel for the carrier gas toflow therethrough in an accelerated state.

According to the gist of the nozzle used in the present invention, adivergent angle of the divergent region formed on the downstream side ofthe throat portion is not larger than 15° in terms of a half-cone angle.

The length of the divergent region is a flight distance untilsolidification or semi-solidification of the atomized metal particlesand is determined on the basis of a flight distance which is determinedby modeling both flight distance of the atomized metal particles and thetemperature of the metal particles. More specifically, the flightdistance until solidification or semi-solidification of the atomizedmetal particles is determined by first determining a flight time untilchange of the atomized metal particles into a solidified orsemi-solidified state and then substituting the said flight time intothe following expression, and the length of the divergent region, in thegist thereof, is set to a length of not shorter than the said flightdistance: $\begin{matrix}{l_{f} = {{u_{g}t_{f}} - {\frac{{u_{g}^{2}\rho_{g}t_{f}} + {\rho_{s}d_{s}a_{g}}}{u_{g}^{2}\rho_{g}}\sqrt{\frac{u_{g}^{2}\rho_{s}d_{s}a_{g}}{{u_{g}^{2}\rho_{g}t_{f}} + {\rho_{s}d_{s}a_{g}}}}} + \frac{\rho_{s}d_{s}a_{g}}{u_{g}\rho_{g}}}} & (18)\end{matrix}$where l_(f) is the flight distance of the particles, t_(f) is the flighttime until solidification or semi-solidification of the particles, u_(g)is flow velocity of gas, p_(g) is gas density, p_(s) is particledensity, d_(s) is particle diameter, and a_(g) is sound velocity of gas.

Preferably, given that an inlet pressure of the carrier gas is p₀ and anozzle outlet pressure thereof is P_(B), the carrier gas is introducedinto the nozzle in a state in which the inlet pressure P₀ satisfies thefollowing expression: $\begin{matrix}{p_{0} \geq {p_{B}\left( {1 + {\frac{\kappa - 1}{2}M^{2}}} \right)}^{\frac{\kappa}{\kappa - 1}}} & (1)\end{matrix}$where κ: specific heat ratio of compressed gas, M: Mach number in theexpanded nozzle portion on the downstream side of the throat portion.

The thermal spraying system according to the present invention, in thegist thereof, comprises the thermal spraying nozzle device constructedas above, a carrier gas supply unit connected to the nozzle through aconduit to introduce the carrier gas under pressure into the nozzle, asealed chamber accommodating the nozzle and a base material forcollision therewith of the ejected particles, and pressure reducingmeans for reducing the internal pressure of the sealed chamber.

The thermal spraying system according to the present invention, in thegist thereof, comprises the thermal spraying nozzle device constructedas above, a molten metal supply unit connected to the storage sectionthrough a connecting pipe to supply molten metal under pressurecontinuously to the molten metal in the storage section, and a basematerial supply unit for continuous supply of the base material.

The present invention is advantageous in that the thermal sprayingmaterial can be supplied constantly and that it is possible to controlthe state of a film or a deposit.

BRIEF DESCRIPTION OF THE DRAWINGS

FIG. 1 is a perspective view showing the construction of a thermalspraying device according to the present invention;

FIGS. 2(a) and 2(b) are explanatory diagrams showing the definition ofan expanded portion of a nozzle.

FIG. 3 is a graph explaining a relation between Mach number and dragcoefficient.

FIG. 4 is a graph showing nozzle lengths according to particlediameters.

FIG. 5 is an explanatory diagram showing a conventional nozzle divergentangle.

FIG. 6 is an explanatory diagram showing a case where a shock wave isgenerated within the nozzle.

FIG. 7 is an explanatory diagram showing a case where a supersonic flowis formed throughout the entire region of the nozzle.

FIG. 8 is a graph showing a typical example of a nozzle shape.

FIG. 9 is a graph showing a nozzle outlet diameter providing anappropriate expansion.

FIG. 10 is a graph showing nozzle length vs. Mach number at a particlediameter of 20 μm and a throat diameter of 25 mm.

FIG. 11 is a graph showing nozzle length vs. gas temperature/velocitydistribution at a particle diameter of 20 μm and a throat diameter of 25mm.

FIG. 12 is a graph showing nozzle length vs. particletemperature/velocity distribution at a particle diameter of 20 μm and athroat diameter of 25 mm.

FIG. 13 is a graph showing nozzle length vs. Mach number at a particlediameter of 20 μm and a throat diameter of 35 mm.

FIG. 14 is a graph showing nozzle length vs. gas temperature/velocitydistribution at a particle diameter of 20 μm and a throat diameter of 35mm.

FIG. 15 is a graph showing nozzle length vs. particletemperature/velocity distribution at a particle diameter of 20 μm and athroat diameter of 35 mm.

FIG. 16 is a graph showing nozzle length vs. Mach number at a particlediameter of 50 μm and a throat diameter of 25 mm.

FIG. 17 is a graph showing nozzle length vs. gas temperature/velocitydistribution at a particle diameter of 50 μm and a throat diameter of 25mm.

FIG. 18 is a graph showing nozzle length vs. particletemperature/velocity distribution at a particle diameter of 50 μm and athroat diameter of 25 mm.

FIG. 19 is a graph showing nozzle length vs. Mach number at a particlediameter of 50 μm and a throat diameter of 35 mm.

FIG. 20 is a graph showing nozzle length vs. gas temperature/velocity ata particle diameter of 50 μm and a throat diameter of 35 mm.

FIG. 21 is a graph showing nozzle length vs. particletemperature/velocity distribution at a particle diameter of 50 μm and athroat diameter of 35 mm.

FIG. 22 is a graph showing nozzle length vs. Mach number at a particlediameter of 100 μm.

FIG. 23 is a graph showing nozzle length vs. gas temperature/velocitydistribution at a particle diameter of 100 μm.

FIG. 24 is a graph showing nozzle length vs. particletemperature/velocity distribution at a particle diameter of 100 μm.

FIG. 25 is an explanatory diagram showing the construction of a thermalspraying system to be applied to a batch process.

FIG. 26 is an explanatory diagram showing the construction of a thermalspraying system to be applied to a continuous molding process.

FIG. 27 is a diagram corresponding to FIG. 1, showing a nozzle of asecond embodiment according to the present invention.

FIG. 28 is a diagram corresponding to FIG. 1, showing a nozzle of athird embodiment according to the present invention.

FIG. 29 is a diagram corresponding to FIG. 1, showing a nozzle of afourth embodiment according to the present invention.

FIG. 30 is a diagram corresponding to FIG. 1, showing a nozzle of afifth embodiment according to the present invention.

FIG. 31 is a diagram corresponding to FIG. 1, showing a nozzle of asixth embodiment according to the present invention.

FIG. 32 is an explanatory diagram showing the construction of aconventional cold spraying system.

BEST MODE FOR CARRYING OUT THE INVENTION

The present invention will be described in detail hereinunder by way ofembodiments thereof illustrated in the drawings.

FIG. 1 illustrates a basic construction of a thermal spraying nozzledevice according to the present invention.

1. Principle of the Thermal Spraying Nozzle Device

The thermal spraying nozzle device shown in the same figure andindicated at 1 supplies molten metal M directly into a supersonic nozzle(hereinafter referred to simply as “nozzle”) 2.

Within the nozzle 2, gas flows at a supersonic velocity, while themolten metal supplied into the nozzle 2 flows at low speed. A shearingforce acts between the two and so does a surface tension of the moltenmetal, whereby the molten metal is atomized downstream of a throatportion 2 a of the nozzle 2.

Atomized metal particles (hereinafter referred to simply as “particles”)are accelerated within the nozzle 2 and are cooled rapidly into asolidified state. Thus, in the thermal spraying nozzle device 1according to the present invention, the throat portion 2 a wherein theatomizing process is performed and a divergent region 2 b wherein aflying/cooling process follows the atomizing process are formedintegrally with each other.

The particles ejected from the nozzle 2 just after solidification comeinto collision with a base material 3 at a velocity of about 450 m/s.The particles generate heat due to deformation caused by the collisionand a portion thereof rise in temperature up to a level of the meltingpoint thereof or higher, whereby the particles adhere to the basematerial 3 (see the impact depositing process in the figure).

The numeral 4 in the figure denotes a storage section for storing themolten metal M, the storage section 4 having a connecting pipe 4 acommunicating with the nozzle 2.

A front end portion of the connecting pipe 4 a is extended as a moltenmetal outlet pipe 4 b toward the center of a cylindrical hole of thethroat portion 2 a and accelerated carrier gas flows over the outerperiphery of the molten metal outlet pipe 4 b.

The principle of collision of solidified particles against the basematerial 3 is the same as in the conventional cold spraying. That is,collided particles undergo a marked plastic deformation and aredepressed like a crater, affording a compact void-free structure withina film (or a deposit layer). Therefore, a molding material thus formedwith the film need not be subjected to HIP (Hot Isostatic Pressing) as apost-treatment, i.e., application of pressure to remove remaining voids.

In case of using nitrogen gas as carrier gas (hereinafter referred tosimply as “gas”) for generating a supersonic gas flow, it is possible toobtain a molding material having a low oxygen content because oxidationdoes not proceed after collision therewith of the particles. Moreover,since the particles are solidified within only 1 ms as a flight timethereof through the nozzle 2, it is possible to prevent the progress ofnitriding.

Besides, molten metal is used as the thermal spraying material and theparticles thereof are brought into collision with the base material 3 ata temperature slightly lower than the solidifying point thereof.Therefore, in comparison with cold spraying, even when the collision isperformed at a low Mach number (e.g., a Mach number of 2 or so), thesurface temperature of the base material 3 reaches a level of not lowerthan the melting point and thus the particles can be adhered positivelyto the base material 3. The Mach number means gas velocity/soundvelocity.

The nozzle 2 has a nozzle length of the expanded portion set at 100 mmor more and is configured so as to operate in a state in which a totalcarrier gas pressure p₀ satisfies the following expression (1):$\begin{matrix}{p_{0} \geq {p_{B}\left( {1 + {\frac{\kappa - 1}{2}M^{2}}} \right)}^{\frac{\kappa}{\kappa - 1}}} & (1)\end{matrix}$where p₀: total carrier gas pressure (throat upstream-side inletpressure), P_(B): throat outlet back pressure, M: Mach number in thethermal spraying material melting section, κ: specific heat ratio of thecarrier gas.

In accordance with the expression (2) the Mach number M is associatedwith a sectional area A* of the throat portion 6 and an intra-nozzleenlarged sectional area A.

The enlarged sectional area includes a conical enlarged portion of agradually increasing diameter from the narrowest portion A* as thethroat portion toward the downstream side, as shown in FIG. 2(a), and anenlarged portion whose diameter suddenly increases from the narrowestportion A* toward the downstream side and then becomes nearly constant,as shown in FIG. 2(b). $\begin{matrix}{\frac{A}{A^{*}} = {\frac{1}{M}\left\lbrack \frac{{\left( {\kappa - 1} \right)M^{2}} + 2}{\kappa + 1} \right\rbrack}^{\frac{\kappa + 1}{2{({\kappa - 1})}}}} & (2)\end{matrix}$

It is known that if compressed gas having a pressure represented by theexpressions (1) and (2) is supplied to a convergent-divergent (Laval)nozzle, there is formed a supersonic flow up to the expanded portion ofthe nozzle. In the narrowest portion the high-speed gas flow becomesMach 1 (about 340 m/s). Molten metal exposed to such a high-speed gasflow is atomized into fine particles. By Hinze (Honze, J. O.,Fundamentals of the Hydrodynamic Machanism of Splitting in DispersionProcesses, AIChEJ., Vol, No.3, 1955, pp.289-295) it has empirically beenmade clear that this is represented by the following expression (3):$\begin{matrix}{\frac{{\rho_{G}\left( {V_{G} - V_{L}} \right)}^{2}D_{L}}{\sigma} \approx 13} & (3)\end{matrix}$where p_(G): gas density, V_(G): gas velocity in nozzle inlet, V_(L):liquid velocity, D_(L): droplet diameter after division, σ: liquidsurface tension.

If reference is made to aluminum alloy as an example of molten metal andnitrogen gas is supplied at a pressure of 0.8 MPa to the nozzle, aparticle diameter of the aluminum alloy after atomization, which isdetermined from the expression (3), is about 20 μm.

The particles after atomization undergo both accelerating and coolingactions by a supersonic gas flow and are eventually ejected from thenozzle 2 while having a supersonic velocity.

The said acceleration and cooling can be estimated by numericalanalysis. More particularly, a mass, momentum and energy conservationexpression as a quasi-one-dimensional compressive fluid conservationtype representation is solved by making an expression (4) simultaneouswith a particles motion equation (6).

2. Method of Numerical Analysis

(1) First, symbols used in a numerical analysis method to be describedlater are shown below.

-   -   A: sectional area of the nozzle    -   CD: particle drag coefficient    -   C_(p): specific heat    -   D: nozzle diameter    -   d: particle diameter    -   f: wall surface friction coefficient    -   g: gravitational acceleration    -   h: specific enthalpy    -   m: mass flow rate    -   Nu: Nusselt number    -   p; gas pressure    -   Pr: Prandt1 number    -   Q: energy per unit time necessary for heating the nozzle    -   R: gas constant    -   Re: Reynolds number    -   T: temperature    -   u: flow velocity    -   x: distance in nozzle flow direction    -   α: Stefan-Boltzmann constant    -   ε: emissivity    -   κ: specific heat ratio    -   λ: thermal conductivity    -   μ: viscosity coefficient    -   p: density

The following are the meanings of subscripts:

-   -   g: gas    -   s: second phase (droplet, particle, powder)    -   x: distance from the nozzle throat portion    -   W: nozzle wall surface        (2) Dominance Equation of Gas Phase

A mass, momentum and energy conservation expression as aquasi-one-dimensional compressive fluid conservation type representationis shown below as an expression (4). $\begin{matrix}{{\frac{\partial U}{\partial t} + \frac{\partial F}{\partial x}} = S} & (4)\end{matrix}$

Provided, however, that the following equation (5) of Johnson-Rubeshinis used in connection with turbulent flow heat transfer of the nozzlewall: $\begin{matrix}{{{U = \begin{pmatrix}{\rho_{g}A} \\{\rho_{g}{Au}_{g}} \\{\rho_{g}{AE}}\end{pmatrix}},{F = \begin{pmatrix}{\rho_{g}{Au}_{g}} \\{{\rho_{g}{Au}_{g}^{2}} + {p\quad A}} \\{\rho_{g}{AU}_{g}H}\end{pmatrix}},{S = \begin{pmatrix}0 \\{{p\quad\frac{\partial A}{\partial x}} - {\pi\quad D\quad f\quad\frac{1}{2}\rho_{g}u_{g}^{2}} + s} \\{{\pi\quad D\quad{Nu}_{x}\frac{\lambda}{x}\left( {T_{W} - T_{g}} \right)} + e}\end{pmatrix}}}{{E = {{\frac{1}{2}u_{g}^{2}} + {\frac{1}{\kappa - 1}\frac{p}{\rho_{g}}}}},{H = {E + \frac{p}{\rho_{g}}}}}{Nu}_{x} = {0.0296\quad\Pr^{\frac{2}{3}}{Re}_{x}^{\frac{4}{5}}}} & (5)\end{matrix}$

In the above expression, s and e stands for a momentum generation termand an energy generation term, respectively, which represent aninteraction between gas phase and second phase.

In actual solution of the expression (1), an advection term is madediscrete using Roe's Flux difference Splitting method which has beenmade into MUSCL (Monotone Upstream-centred Schemes for Conversion laws),and time integral is performed using a four-stage Runge-Kutta method.

(3) Dominance Equation of the Second Phase (Droplet, Particle, Powder)

Particles' velocity can be obtained by solving the following particlemotion equation (6): $\begin{matrix}{{{\frac{\partial u_{s}}{\partial t} + {u_{s}\frac{\partial u_{s}}{\partial x}}} = {{\frac{\rho_{s} - \rho_{g}}{\rho_{s}}g} - {\frac{u_{s}}{{\overset{.}{m}}_{s}}s}}}{{Provided},{{however}\text{:}}}} & (6) \\{s = {\frac{3}{2}\frac{{\overset{.}{m}}_{s}C_{D}}{d_{s}\rho_{s}u_{s}}\frac{1}{2}{\rho_{g}\left( {u_{s} - u_{g}} \right)}{{u_{s} - u_{g}}}}} & (7)\end{matrix}$

The following equation (8) of Kurten is used for drag coefficient:$\begin{matrix}{{C_{D} = {0.28 + {6\quad{Re}^{- 0.5}} + {21\quad{Re}^{- 1}}}}{{Re} = \frac{\rho_{g}{{u_{s} - u_{g}}}d_{s}}{\mu}}} & (8)\end{matrix}$

The particle temperature can be obtained by solving the followingparticle energy equation (9): $\begin{matrix}{{\frac{\partial h_{s}}{\partial t} + {u_{s}\frac{\partial h_{s}}{\partial x}}} = {{- \frac{u_{s}}{{\overset{.}{m}}_{s}}}\left( {q + e} \right)}} & (9)\end{matrix}$

However, in the case of a heat insulating wall with nozzle walltemperature equal to the gas temperature: $\begin{matrix}{{e = {\frac{6\quad{\overset{.}{m}}_{s}}{\rho_{s}u_{s}d_{s}}\left\{ {{{Nu}\quad\frac{\lambda}{d_{s}}\left( {T_{s} - T_{g}} \right)} + {\alpha\quad{ɛ\left( {T_{s}^{4} - T_{W}^{4}} \right)}}} \right\}}},{q = 0}} & (10)\end{matrix}$

In case of the nozzle wall 1 b being a heated isothermal wall:$\begin{matrix}{{e = {\frac{6\quad{\overset{.}{m}}_{s}}{\rho_{s}u_{s}d_{s}}{Nu}\quad\frac{\lambda}{d_{s}}\left( {T_{s} - T_{g}} \right)}},{q = {\frac{6\quad{\overset{.}{m}}_{s}}{\rho_{s}u_{s}d_{s}}\alpha\quad{ɛ\left( {T_{s}^{4} - T_{W}^{4}} \right)}}}} & (11)\end{matrix}$

The following expression (12) of Ranz-Marshall is used for Nusseltnumber: $\begin{matrix}{{Nu} = {2 + {0.6\quad\Pr^{\frac{1}{3}}{Re}^{\frac{1}{2}}}}} & (12)\end{matrix}$

In actual solution of the expressions (6) and (9), QUICK method is usedfor making the advection term discrete and time integral is performedusing a four-stage Runge-Kutta method.

(4) Quantity of Heat Required for Heating the Nozzle

The quantity of heat necessary for maintaining the isothermal conditioncan be estimated by the following expression (13): $\begin{matrix}{Q = {\int_{0}^{L}{\left\lbrack {{\pi\quad D\quad{Nu}_{x}\frac{\lambda}{x}\left( {T_{W} - T_{g}} \right)} - q} \right\rbrack{\mathbb{d}x}}}} & (13)\end{matrix}$(5) Nozzle Length

In operation using the thermal spraying nozzle device according to thepresent invention, the distance from the nozzle outlet to the deposit isset extremely short because the device is configured so that theatomized particles collide with the deposit before deceleration of theparticles' velocity. Therefore, it is approximately presumed that theparticle velocity and enthalpy in the nozzle output are substantiallymaintained, allowing the particles to be deposited.

The state of the deposit depends much on the state of the particlesbeing deposited, but in case of the particles being brought intocollision and deposited at a subsonic velocity as in the conventionalthermal spraying nozzle device, the particles cannot be adhered to thebase material or the deposit if the particles are in a solidified state.

On the other hand, the thermal spraying nozzle device according to thepresent invention defines as an operating condition that semi-solidifiedor solidified particles with a greater solid phase ratio so far notutilized should be brought into collision and deposited at a supersonicvelocity. In this connection, it is presumed that molten metal changesinto a semi-solidified state while being atomized and flying, and aminimum flight distance required until that time-point is determined.This flight distance is presumed to be a minimum nozzle length requiredfor the device.

As noted earlier, a motion equation which represents acceleration of theparticles is the following equation (6): $\begin{matrix}{{\frac{\partial u_{s}}{\partial t} + {u_{s}\frac{\partial u_{s}}{\partial x}}} = {{\frac{\rho_{s} - \rho_{g}}{\rho_{s}}g} - {\frac{u_{s}}{{\overset{.}{m}}_{s}}s}}} & (6)\end{matrix}$

The expression (6) is convenient for numerical calculation using a fixedcalculation lattice because it is described from the Euler's coordinatesystem which stands still together with the nozzle.

However, the expression in question is inconvenient for tracing thestate of the particles one by one to check both particle velocity andparticle enthalpy. Therefore, if it is expressed in terms of an equationdescribed from the Lagrangian coordinate system which moves togetherwith flying particles, the following equation (14) is obtained.

However, for the purpose of simplification, a gravitational term whichexerts little influence is ignored. Further, it is presumed that in asection where the flight distance is still short the particles are inthe course of acceleration and are pushed from behind like a fair windby the gas flow and that there always exists the relationship of gasflow velocity u_(g)> particle velocity u_(s). $\begin{matrix}{\frac{\mathbb{d}u_{s}}{\mathbb{d}t} = {\frac{3}{2}\frac{C_{D}}{d_{s}\rho_{s}}\frac{1}{2}{\rho_{g}\left( {u_{g} - u_{s}} \right)}^{2}}} & (14)\end{matrix}$

By taking a relative velocity between the gas flow velocity u_(g) andthe particle velocity us to obtain U=u_(g)−u_(s) and by assuming thatthe gas flow velocity as a supersonic velocity is approximately constantwithin the nozzle, the expression (14) can be transformed into thefollowing expression (15): $\begin{matrix}{\frac{\mathbb{d}U}{\mathbb{d}t} = {{- \frac{3}{2}}\frac{C_{D}}{d_{s}\rho_{s}}\frac{1}{2}\rho_{g}U^{2}}} & (15)\end{matrix}$

In the expression (15), the drag coefficient C_(D) can be expressed by afunction of Reynolds number in the case where the relative velocity U isa subsonic velocity, as shown in the expression (12), but in the statejust after atomization the relative velocity U is very likely to be asupersonic velocity. Therefore, in the graph of FIG. 3 (an explanatorydiagram of a sphere obtained from bullet route measurement, as well asdrag coefficient of cone-cylinder and Mach number dependence), the dragcoefficient in question is approximated by the following expression (16)from an experimental result on drag coefficient of Mach number and thesphere (see an approximate line E in the figure).

FIG. 3 was quoted from “2nd edition McGraw-Hill Series in Aeronauticaland Aerospace Engineering, Modern Compressible Flow with historicalPerspective.” $\begin{matrix}{C_{D} = {{\frac{2}{3}M} = {\frac{2}{3}\frac{U}{a_{g}}}}} & (16)\end{matrix}$where a_(g) stands for the sound velocity of gas and M stands for Machnumber.

From the expressions (16) and (15) there is obtained the followingexpression (17) which expresses a relation between the particle flighttime, t, and the relative velocity U: $\begin{matrix}{U = \sqrt{\frac{u_{g}^{2}\rho_{s}d_{s}a_{g}}{{u_{g}^{2}\rho_{g}t} + {\rho_{s}d_{s}a_{g}}}}} & (17)\end{matrix}$where the particle velocity us is assumed equal to zero at t=0.

The relation between the flight time t_(f) and the flight distance l_(f)is obtained from the following expression (18): $\begin{matrix}{l_{f} = {{u_{g}t_{f}} - {\frac{{u_{g}^{2}\rho_{g}t_{f}} + {\rho_{s}d_{s}a_{g}}}{u_{g}^{2}\rho_{g}}\sqrt{\frac{u_{g}^{2}\rho_{s}d_{s}a_{g}}{{u_{g}^{2}\rho_{g}t_{f}} + {\rho_{s}d_{s}a_{g}}}}} + \frac{\rho_{s}d_{s}a_{g}}{u_{g}\rho_{g}}}} & (18)\end{matrix}$where u_(g) is the flow velocity of gas, p_(g) is the density of gas,p_(s) is the density of particles, and d_(s) is particle diameter.

Once the flight time t_(f) until the particles reach a semi-solidifiedstate is known, the particle flight distance l_(f) in the past iscalculated, corresponding to the required minimum nozzle length. Forthis reason, the flight time t_(f) until the particles becomesemi-solidified is determined. Cooling of the particles is given interms of the foregoing expression (9): $\begin{matrix}{{\frac{\partial h_{s}}{\partial t} + {u_{s}\frac{\partial h_{s}}{\partial x}}} = {{- \frac{u_{s}}{{\overset{.}{m}}_{s}}}\left( {q + e} \right)}} & (9)\end{matrix}$

Like the expression (14), this can be expressed in terms of theLagrangian coordinate system as follows (19): $\begin{matrix}{\frac{\mathbb{d}h_{s}}{\mathbb{d}t} = {\frac{6}{\rho_{s}d_{s}}\left\{ {{{Nu}\quad\frac{\lambda}{d_{s}}\left( {T_{g} - T_{s}} \right)} + {\alpha\quad{ɛ\left( {T_{W}^{4} - T_{s}^{4}} \right)}}} \right\}}} & (19)\end{matrix}$

Initial molten metal temperature, liquidus temperature and solidustemperature are almost equal approximately and if the value there of isrepresented by the material melting point T_(m), T_(s)=T_(m). The gastemperature T_(s) and the nozzle wall surface temperature T_(w) are alsoassumed equal approximately.

Nusselt number Nu, which represents the degree of heat transfer, isrepresented by the expression (12), but can be rewritten into thefollowing expression (20) using the relative velocity U: $\begin{matrix}{{Nu} = {2 + {0.6\quad{\Pr^{\frac{1}{3}}\left( \frac{\rho_{g}d_{s}}{\mu_{g}} \right)}^{\frac{1}{2}}U^{\frac{1}{2}}}}} & (20)\end{matrix}$

If latent heat of solidification of the molten metal is assumed to be L,the following expression (21) is valid in order to attain asemi-solidified state with a larger solid phase ratio: $\begin{matrix}{{\int_{0}^{t_{1}}{{- \frac{\mathbb{d}h}{\mathbb{d}t}}\quad{\mathbb{d}t}}} \geq \frac{L}{2}} & (21)\end{matrix}$

In the above expression there is written L/2 because a nearlyintermediate point of transition from liquid phase to solid phasecorresponds to a semi-solidified state.

In this case there is established an equal sign in the expression (21)because determining a minimum nozzle length means determining theshortest flight time t_(f) until the particles becomes semi-solidified.

If Nusselt number Nu is eliminated from the expressions (19) and (20)and the relative velocity U is also eliminated using the expression (17)and if an equal sign expression in the expression (21) is used, there isobtained the following relationship (22) of the shortest flight timet_(f) until the particles reaches a semi-solidified state:$\begin{matrix}{{0.6\quad\Pr^{\frac{1}{3}}{\lambda\left( {T_{m} - T_{g}} \right)}}{{{\sqrt{\frac{\rho_{g}}{\mu_{g}d_{s}}}\begin{Bmatrix}{{\frac{4\left( {{u_{g}^{2}\rho_{g}t_{f}} + {\rho_{s}d_{s}a_{g}}} \right)}{3u_{g}^{2}\rho_{g}}\left( \frac{u_{g}^{2}\rho_{s}d_{s}a_{g}}{{u_{g}^{2}\rho_{g}t_{f}} + {\rho_{s}d_{s}a_{g}}} \right)^{0.25}} -} \\\frac{4\rho_{s}d_{s}a_{g}}{3u_{g}^{\frac{3}{2}}\rho_{g}}\end{Bmatrix}} + {\left\{ {{\frac{2\lambda}{d_{s}}\left( {T_{m} - T_{g}} \right)} + {\alpha\quad{ɛ\left( {T_{m}^{4} - T_{W}^{4}} \right)}}} \right\} t_{f}}} = \frac{\rho_{s}d_{s}L}{12}}} & (22)\end{matrix}$where Pr is Prandt1 number of gas, λ is the thermal conductivity of gas,T_(m) is the material melting point, T_(g) is the temperature of gas,and μ_(g) is the viscosity coefficient of gas.

The above expression (22) cannot be solved for t_(f), but can be solvednumerically using the Newton's method.

Thus, by determining the shortest flight time t_(f) from the expression(22) and substituting it into the expression (18) there is determinedthe shortest flight time, i.e., minimum nozzle length l_(f).

The thermal spraying nozzle device according to the present invention ischaracterized by being a device using a nozzle with a length of notsmaller than the above nozzle length l_(f), and by accelerating theparticles up to a supersonic velocity the particles even in a solidifiedstate adhere to the base material or deposit. Thus, the nozzle lengthhas no upper limit theoretically.

FIG. 4 is a graph of having determined minimum nozzle lengths concretelywith use of aluminum and copper. The nozzle lengths shown therein areconsidered necessary when particles of various diameters assume asemi-solidified state with a solid phase ratio exceeding 0.5. In thesame graph, the particle diameter is plotted along the axis of abscissaand the nozzle length along the axis of ordinate. Carrier gas conditionsare the same as in Table 1 which will be described later.

As a result of atomization, when an average particle diameter in termsof volume occupancy for example is 50 μm, a required nozzle length is0.17 m in case of aluminum and 0.12 m in case of copper.

Heretofore, when flowing molten metal into a supersonic nozzle for thepurpose of atomization, there is used a nozzle having a larger divergentangle (a divergent angle of the divergent region on the downstream sideof the throat portion) of θ>15° in terms of a half-cone angle as shownin FIG. 5 in order to avoid adhesion of the particles to the inner wallsurface of the nozzle. The said half-cone angle means the angle betweenthe central nozzle axis and the nozzle inner wall.

In this case, the sectional area ratio A/A* increases abruptly and sodoes Mach number (see the expression (2)), but a shock wave frontappears upon arrival at Mach number M₁ which is determined from theisentropic change expression (23) and the vertical shock waverelationship (24), and with this as a boundary the gas flow on thedownstream side becomes a subsonic flow and the divergent angle of thenozzle inner wall is large, so that the gas flow near the inner wallsurface peels off the inner wall surface.

At this time, the Mach number M₁ is determined from the expression (25)and the sectional area ratio A/A* at the position where the shock wavefront appears is determined from the expression (26).

Such a nozzle has heretofore been suitable for atomization, but theintra-nozzle gas flow immediately becomes a subsonic flow and theconcept of accelerating particles is not existent. On the other hand, inthe construction of the nozzle according to the present invention, theparticles after atomization are accelerated up to a supersonic velocitywhile setting the divergent angle of the nozzle at 15° or less toprevent separation of the gas flow and so that the particles even in asemi-solidified state can be adhered to the base material or deposit. Inother words, in the nozzle according to the present invention, thedistance from the narrowest portion of the nozzle up to the shock wavefront generating position is extended long until the particles reach asolidified or semi-solidified state. $\begin{matrix}{\frac{p_{0}}{p_{1}} = \left( {1 + {\frac{\kappa - 1}{2}M_{1}^{2}}} \right)^{\frac{\kappa}{\kappa - 1}}} & (23) \\{\frac{p_{1}}{p_{B}} = \frac{{2\quad\kappa\quad M_{1}^{2}} - \left( {\kappa - 1} \right)}{\kappa + 1}} & (24) \\{\frac{p_{0}}{p_{B}} = {\frac{{2\quad\kappa\quad M_{1}^{2}} - \left( {\kappa - 1} \right)}{\kappa + 1}\left( {1 + {\frac{\kappa - 1}{2}M_{1}^{2}}} \right)^{- \frac{\kappa}{\kappa - 1}}}} & (25) \\{\frac{A_{1}}{A^{*}} = {\frac{1}{M_{1}}\left\lbrack \frac{{\left( {\kappa - 1} \right)M_{1}^{2}} + 2}{\kappa + 1} \right\rbrack}^{\frac{\kappa + 1}{2{({\kappa - 1})}}}} & (26)\end{matrix}$

In accordance with the above description, conditions for the supersonicnozzle in the present invention can be defined by the following (a) to(c):

(a) The divergent angle of the nozzle should be θ≦15° in terms of ahalf-cone angle.

(b) The divergent angle of the nozzle should be θ≦15° in terms of ahalf-cone angle, and when a shock wave upstream Mach number M₁ isdetermined by the expression (25) on the basis of the total carrier gaspressure p₀ and the nozzle outlet back pressure P_(B) and is substitutedinto the expression (26) to determine the sectional area A₁ of thenozzle, the nozzle length l_(f) up to the position corresponding to thesectional area A₁ of the nozzle should be not shorter than a minimumnozzle length l_(f) determined from both the expression (18) and therelationship (22) which defines the shortest flight time until theparticles become semi-solidified.

FIG. 6 shows a case where a shock wave is generated within the nozzle.

(c) The divergent angle of the nozzle should be θ≦15° in terms of ahalf-cone angle, the nozzle length l_(f) should be not shorter than theshortest nozzle length l_(f) determined from both the expression (18)and the relationship (22) which defines the shortest flight time untilthe particles become semi-solidified, and when a shock wave upstreamMach number M₁ is determined by the expression (25) on the basis of thetotal carrier gas pressure p₀ and the nozzle outlet back pressure P_(B)and is substituted into the expression (26) to determine the sectionalarea A₁ of the nozzle, the sectional area A₁ should be larger than thenozzle outlet sectional area A_(e).

In this case, since a supersonic flow is generated in the whole regionof the nozzle as shown in FIG. 7, a shock wave front is generated on thedownstream side of the nozzle outlet.

3. Designing an Actual Nozzle

3-1) Physical Property Values of Materials and Restraint Conditions

Physical property values of materials and restraint conditions usedcalculating an actual nozzle are shown in Table 1. TABLE 1 Physicalproperty values of materials and restraint conditions used incalculating an actual nozzle Material/ Physical Properties Kind Shape &Conditions Value Unit Carrier Nitrogen Specific heat 297 J/KgK GasSpecific heat ratio 1.4 Thermal conductivity 2.5 × 10⁻³ W/mK Viscositycoefficient  18 × 10⁻⁶ Pas Prandtl number 0.72 Initial total 293 Ktemperature Total pressure 0.8 Mpa Back pressure 0.1 MPa ParticlesAluminum Density 2700 kg/m³ alloy emissivity 0.5 Specific heat in 902J/kgK liquid phase Specific heat in 951 J/kgK solid phase Latent heat of398 × 10³  J/Kg solidification Liquefaction start 934 K pointtemperature Solidification start 773 K point temperature Initialtemperature 1173 K Initial flow velocity 6 m/s Nozzle Axisym- Maximum 5deg metric half-cone angle specific heat

By the maximum half-cone angle in the above table is meant a maximumangle between the nozzle axis and the nozzle inner wall.

3-2) Conditions for Study

-   -   Molten metal (particles) mass flow rate [kg/s],    -   four conditions: 0.025, 0.050, 0.075, 0.100    -   Particle diameter [μm], three conditions: 20, 50, 100    -   Nozzle throat portion dia. [mm], two conditions: 25, 35

The mass flow rate of gas at the throat diameter of 25 mm and that atthe throat diameter of 35 mm correspond to 0.9 [kg/s] and 1.8 [kg/s],respectively.

Under the above conditions, a nozzle shape in case of an appropriateexpansion (static pressure in nozzle outlet=back pressure=atmosphericpressure) being obtained was determined and the relation betweenparticle temperature and particle velocity was checked. At a supersonicflow there is no influence exerted on the upstream side from thedownstream side and therefore, for example, the result of calculation atthe position of 300 mm in a nozzle 500 mm long can be regarded as it isas the state in the outlet of a 300 mm long nozzle. This is a differentpoint from a subsonic nozzle.

3-3) Construction of the Actual Nozzle

3-3-1) Entire Construction

A typical example of a nozzle shape designed for spray acceleration isshown in the graph of FIG. 8.

In the illustrated example the maximum half-cone angle of the nozzle isset at 5° (see Table 1).

This nozzle is configured with a view to (a) expanding disperseddroplets after atomization quickly up to the maximum diameter so as notto adhere to the nozzle wall and (b) taking long a straight pipe portionat the maximum diameter at which the velocity becomes maximum so as toaccelerate the particles.

However, in comparison with a conical nozzle used commonly in coldspray, the nozzle of this embodiment is inconvenient in that the wholeof a straight pipe portion which occupies most of the nozzle becomessubsonic in the case where a pressure ratio is lower than a design valueor when a lot of cold particles are supplied. Thus, the nozzle of thisembodiment is unsuitable for operation in a deviated state from thedesign value, but is suitable for production equipment in whichoperation is repeated under the same conditions. In this connection, thegraph of FIG. 9 shows nozzle outlet diameters affording an appropriateexpansion on the premise that operation is performed under the sameconditions as just referred to above.

In the same graph, the reason why the nozzle outlet diameter increaseswith an increase in flow rate of molten metal no matter which of 25 mmand 35 mm of the nozzle throat diameter may be is that the gas receivesthe heat carried in by the molten metal, creating an expandable state.

An interesting point is that if a nozzle is designed under the conditionof a small mass flow rate of molten metal, even if the molten metal issupplied to the nozzle at a flow rate exceeding the design value,operation can be done up to a limitation based on a momentum deliveredto the particles although the acceleration efficiency decreases due todeficient expansion. Conversely, it is seen that at a mass flow rate ofmolten material smaller than value it is impossible to effectacceleration up to a supersonic velocity.

Next, Table 2 shows a relation between nozzle throat diameters resultingfrom design calculation in the actual nozzle and mass flow rate of gasin case of heating being not performed. TABLE 2 Results of nozzle designcalculation of this time and mass flow rate of gas Nozzle Mass Flow Rateof Throat Outlet Mass Flow Rate of Gas, kg/s Molten Metal Dia. Dia.Particle Dia., μm kg/s mm mm 20 50 100 0 25 32 0.91 35 45 1.79 0.025 2534 0.91 0.91 35 47 1.79 1.79 0.05 25 36 0.90 0.91 0.92 35 48 1.78 1.790.075 25 38 0.88 Subsonic 35 49 1.76 1.79 1.79 0.1 25 Decelerate tosubsonic immediately influx 35 50 1.76 1.79 1.793-3-2) In Case of the Particle Diameter after Atomization being 20 μm:

In FIGS. 10 to 12 there are shown intra-nozzle Mach numberdistributions, gas temperature/velocity distributions, and particletemperature/velocity distributions, respectively, assuming that theparticle diameter after atomization is 20 μm and the nozzle throatdiameter is 25 mm. In the graphs to be described below, Distance plottedalong the axis of abscissa represents the nozzle length, while in theaxis of ordinate, Mach number, Gas temp, Gas Velc, Solid temp, and SolidVelc, represent Mach number, gas temperature, gas velocity, particletemperature, and particle velocity, respectively.

In FIGS. 13 to 15 there are shown intra-nozzle Mach numberdistributions, gas temperature/velocity distributions, and particletemperature/velocity distributions, respectively, assuming that theparticle diameter after atomization is 20 μm and the nozzle throatdiameter is 35 mm.

Because of a heated Rayleigh flow which receives heat from molten metal,Mach number decreases, gas temperature rises, and gas velocitydecreases.

Since in this embodiment the nozzle outlet diameter is determined so asto give an appropriate expansion after heating, the static pressure ofgas is almost equal to the atmospheric pressure and gas velocities areall 510 m/s or so.

It is interesting to note that if the nozzle outlet diameter isdetermined so as to give an appropriate expansion after heating withrespect to each of such conditions, the state on the particles side nowaffords almost equal results in both particle velocity and particletemperature.

This is because intra-nozzle gas velocity distributions are almost equaland the gas temperature difference is small in comparison with thedifference in temperature from molten metal.

The difference between the throat diameters 25 mm and 35 mm appears inthe gas temperature, but does not appear in the gas velocity, as shownin FIGS. 11 and 14. Therefore, as to the particles influenced by the gastemperature, a difference appears in the particle temperature, but doesnot appear in the particle velocity.

In the case where the particle diameter is 20 μm, solidification iscompleted at a nozzle length of about 160 mm, but the particle velocityis only about 400 m/s. In this case, if the nozzle length is extended to500 mm, it is possible to accelerate the particle velocity to 480 m/s,but the particles are cooled to a temperature of 400K.

Thus, in case of the particle diameter being 20 μm, there is a tendencythat the particles are cooled too much in comparison with acceleration,so it is necessary to determine the nozzle length prudently.

3-3-3) In Case of the Particle Diameter after Atomization being 50 μm:

In FIGS. 16 to 18 there are shown intra-nozzle Mach numberdistributions, gas temperature/velocity distributions, and particletemperature/velocity distributions, respectively, assuming that theparticle diameter after atomization is 50 μm and the nozzle throatdiameter is 25 mm.

Likewise, in FIGS. 19 to 21 there are shown intra-nozzle Mach numberdistributions, gas temperature/velocity distributions, and particletemperature/velocity distributions, respectively, assuming that theparticle diameter after atomization is 50 μm and the nozzle throatdiameter is 35 mm.

Mach number, gas temperature and particle velocity show a tendency notgreatly different from that in the case of the particle diameter being20 μm, but a conclusive different point resides in the particletemperature cooling velocity shown in FIGS. 18 and 21.

In the case where the particle diameter is 50 μm, a flight distance ofabout 1.2 m is needed within the nozzle until completion ofsolidification. If the nozzle length is extended to 1.2 m accordingly,an asymptotic line of particle acceleration is fairly approachedconveniently.

In this condition the particles are ejected from the nozzle at aparticle temperature of 750K and a particle velocity of 470 m/s and thusthis condition is most preferred as an impact depositing condition forthe base material.

3-3-4) In Case of the Particle Diameter after Atomization being 100 μm:

In FIGS. 22 to 24 there are shown intra-nozzle Mach numberdistributions, gas temperature/velocity distributions, and particletemperature/velocity distributions, respectively, in case of theparticle diameter after atomization being 100 μm.

This calculation result shows that a nozzle length of 5 m is neededuntil solidification after a lowering of the cooling velocity in case ofthe particle diameter being 100 μm. Since particle acceleration hasalready ended at the time point corresponding to the nozzle length of 3m and the particle velocity now reaches about 450 m/s, cooling becomeslater. Such a situation occurs when atomization cannot be done to asatisfactory extent.

FIG. 25 shows a construction in case of a thermal spraying systemaccording to the present invention being applied to a batch process.

In the same figure, the same constituent elements as in FIG. 1 areidentified by the same reference numerals, and explanations thereof willbe omitted.

As carrier gas, helium gas of a low molecular weight, which is preferredin point of a sound velocity becoming high when accelerating particles,is used in place of nitrogen gas.

Carrier gas supplied from a helium gas cylinder 10 is branched into twopipes 11 and 12. The carrier gas flowing in the pipe 11 imparts a headpressure to molten metal stored in a storage section 4, while thecarrier gas flowing in the pipe 12 is introduced into a nozzle 2 andpasses through a throat portion 2 a of the nozzle 2, whereby it isaccelerated to a supersonic velocity. The helium gas cylinder 10 and thepipes 11, 12 function as a carrier gas supply unit for the supply ofcarrier gas under pressure.

The molten metal flowing down from the storage section 4 is atomized bythe supersonic gas flow in the nozzle 2, then the atomized particles arecooled in the nozzle 2 and ejected from a front end of the nozzle 2.

The ejected particles collide with and adhere to the surface of a basematerial 3. The nozzle 2 and the base material 3 are accommodated withina chamber 13 which is a sealed chamber. The chamber 13 is connected to agas storage tank 16 via a cyclone unit 14 as an exhaust unit and anexhaust vacuum pump (pressure reducing means) 15. The cyclone unit 14recovers particles suspended in exhaust air and supplies only gas to theexhaust vacuum pump 15.

The exhaust unit is provided for increasing the Mach number of carriergas and thereby increasing the particle velocity. The helium gasrecovered into the gas storage tank 16 is compressed by a compressor 17and is re-utilized.

FIG. 26 shows a basic construction in case of applying a thermalspraying system according to the present invention to a continuousmolding process.

In the continuous molding process shown in the same figure, a continuousmelting furnace 20 is connected to a storage section 4 and the storagesection 4 and the continuous melting furnace 20 are in communicationwith each other through a connecting pipe 21. The height of thecontinuous melting furnace 20 is set so that the inner pressure of thestorage section 4 is held at 0.8 MPa by a head pressure. The continuousmelting furnace 20 disposed at the above predetermined height functionsas a molten metal supply unit for continuous supply of molten metalunder pressure.

Thus, molten metal can be supplied to a nozzle 2 continuously from thestorage section 4.

While a base material 22 rotates also in the direction of arrow A, it isdrawn out in the direction of arrow B by rotation of take-off rollers(base material supply unit) 23 a and 23 b, whereby particles can besprayed continuously onto the base material 22 to effect molding.

FIGS. 27 to 31 show other embodiments of nozzles 2 according to thepresent invention. In each of the nozzles, the nozzle itself isfabricated using a non-metal such as a ceramic material or carbon todeteriorate the surface affinity, whereby metal particles adhered to theinner surface of the nozzle can be blown off easily by a supersonic gasflow. In those figures, the same constituent elements as in FIG. 1 areidentified by the same reference numerals, and explanations thereof willbe omitted.

In the case of a nozzle 40 shown in FIG. 27, a nozzle 41 is fabricatedusing zirconium for thermal spraying of aluminum alloy, the outside ofthe nozzle 41 being covered with a ceramic cylinder 42, and a nozzleheater 43 capable of raising temperature up to a maximum of 900° C. iswound plural turns round the cylinder 42. As the material of the nozzle41 it is preferable to use a material called partially stabilizedzirconium with for example yttria (Y₂O₃) added as a stabilizing agent,which material possesses high strength, high abrasion resistance andhigh corrosion resistance.

As to a nozzle 44 shown in FIG. 28, the nozzle itself is constituted bya ceramic fiber heater 45. More specifically, a material consistingprincipally of alumina and silica is made into a high temperatureinsulating ceramic fiber, followed by embedding a heating element intothe ceramic fiber and subsequent integral molding. Numerals 46 a and 46b in the figure denote electrode connecting portions of the heater.

According to the construction of a nozzle 47 shown in FIG. 29, a carbonheater 49 is disposed around an outer wall of a body portion of aceramic nozzle 48 and heating is performed by radiation.

The carbon heater 49 is divided into plural portions by slits 51 d and51 e which are formed a predetermined length alternately from both upperand lower sides of a cylindrical nozzle 48. Numerals 49 a and 49 bdenote electrode connecting portions of the carbon heater 49. Numeral 50denotes a cylindrical reflection case having a specular-finished innerwall and it is provided for enhancing the radiation efficiency.

In the nozzle 47 constructed as above, when electric power is fed from apower supply (not shown) to the carbon heater 49 via the electrodeconnecting portions 49 a and 49 b, the carbon heater 49 generates heatfrom the interior thereof due to Joule heat induced by the supply ofelectric power. As a result, the ceramic nozzle 48 is heated byradiation heat transfer from the carbon heater 49 and the metal adheredto the inner wall of the nozzle 37 is melted.

As to a nozzle 51 shown in FIG. 30, the nozzle itself is fabricated by acarbon heater 52. Numerals 52 a and 52 b denote electrode connectingportions of the carbon heater. By replacing a ceramic nozzle with acarbon or carbon composite nozzle, the radiation rate of the nozzlesurface is further enhanced and it is possible to further improve theheating efficiency of the nozzle 51.

In FIGS. 29 and 30, the presence of oxygen causes an oxidation reactionof carbon itself, so for avoiding such an inconvenience, the whole ofthe system is covered with a chamber and gas such as argon or nitrogengas is used as high pressure gas to purge the interior of the chamberwith an inert atmosphere.

Also by fabricating a nozzle with a metallic material superior inthermal conductivity, e.g., copper, and thermal spraying of a ceramicmaterial to the inner wall of the nozzle thus fabricated to form aceramic film, it is possible to deteriorate the surface affinity as isthe case with each of the foregoing nozzles.

In the case of a nozzle 53 shown in FIG. 31, a zirconium film (theportion indicated by a thick broken line in the figure) 55 is formed onan inner surface of a copper nozzle 54 and a nozzle heater 43 is woundplural turns round an outer periphery surface of the nozzle.

INDUSTRIAL APPLICABILITY

The thermal spraying nozzle device and the thermal spraying systemaccording to the present invention are suitable in a field in which itis required to supply a thermal spraying material constantly onto a basematerial and control the state of a film or deposit formed on the basematerial.

1. A thermal spraying nozzle device wherein carrier gas is introducedfrom an inlet side of a nozzle to form a supersonic gas flow and athermal spraying material is atomized and ejected by said gas flow, saidthermal spraying nozzle device comprising, a storage section storingmolten metal as said thermal spraying material connected to an end onthe inlet side of said nozzle through a connecting pipe, and, saidnozzle having a throat portion and a divergent region in a downstream ofsaid throat portion toward an outlet side to form the supersonic gasflow, wherein said thermal spraying nozzle device is configured suchthat metal particles atomized by the supersonic gas flow are cooled to asolidified or semi-solidified state in said divergent region and thenejecting in a predetermined direction from the outlet side of saidnozzle.
 2. The thermal spraying nozzle device according to claim 1,wherein, within said connecting pipe, a molten metal outlet pipe isextended from said storage section toward the center in said throatportion or the center on the downstream side of the throat portion andan outside portion of said molten metal outlet pipe constitutes achannel for the carrier gas to flow therethrough in an acceleratedstate.
 3. The thermal spraying nozzle device according to claim 1,wherein a divergent angle of said divergent region formed on thedownstream side of said throat portion is not larger than 15° in termsof a half-cone angle.
 4. The thermal spraying nozzle device according toclaim 3, wherein the length of said divergent region is a flightdistance until solidification or semi-solidification of the atomizedmetal particles and is determined on the basis of a flight distancewhich is determined modeling both flight distance of the atomized metalparticles and the temperature of the metal particles.
 5. A thermalspraying nozzle device according to claim 4, wherein the flight distanceuntil solidification or semi-solidification of said atomized metalparticles is determined by first determining a flight time until changeof the atomized metal particles into a solidified or semi-solidifiedstate and then substituting said flight time into the followingexpression, and the length of said divergent region is set to a lengthof not shorter than said flight distance: $\begin{matrix}{l_{f} = {{u_{g}t_{f}} - {\frac{{u_{g}^{2}\rho_{g}t_{f}} + {\rho_{s}d_{s}a_{g}}}{u_{g}^{2}\rho_{g}}\sqrt{\frac{u_{g}^{2}\rho_{s}d_{s}a_{g}}{{u_{g}^{2}\rho_{g}t_{f}} + {\rho_{s}d_{s}a_{g}}}}} + \frac{\rho_{s}d_{s}a_{g}}{u_{g}\rho_{g}}}} & (18)\end{matrix}$ where l_(f) is the flight distance of the particles, t_(f)is the flight time until solidification or semi-solidification of theparticles, u_(g) is flow velocity of gas, p_(g) is gas density, p_(s) isparticle density, d_(s) is particle diameter, and a_(g) is soundvelocity of gas.
 6. The thermal spraying nozzle device according toclaim 1, wherein, given that an inlet pressure of the carrier gas is p₀and a nozzle outlet pressure thereof is P_(B), the carrier gas isintroduced into said nozzle in a state in which the inlet pressure p₀satisfies the following expression: $\begin{matrix}{p_{0} \geq {p_{B}\left( {1 + {\frac{\kappa - 1}{2}M^{2}}} \right)}^{\frac{\kappa}{\kappa - 1}}} & (1)\end{matrix}$ where κ: specific heat ratio of compressed gas, M: Machnumber in the expanded nozzle portion on the downstream side of thethroat portion.
 7. A thermal spraying system comprising: a thermalspraying nozzle device wherein carrier gas is introduced from an inletside of a nozzle to form a supersonic gas flow and a thermal sprayingmaterial is atomized and ejected by said gas flow, said thermal sprayingnozzle device comprising, a storage section storing molten metal as saidthermal spraying material connected to an end on the inlet side of saidnozzle through a connecting pipe, and, said nozzle having a throatportion and a divergent region in a downstream of said throat portiontoward an outlet side to form the supersonic gas flow, wherein saidthermal spraying nozzle device is configured such that metal particlesatomized by the supersonic gas flow are cooled to a solidified orsemi-solidified state in said divergent region and then ejecting in apredetermined direction from the outlet side of said nozzle, a carriergas supply unit connected to said nozzle through a conduit to introducethe carrier gas under pressure into the nozzle; a sealed chamberaccommodating said nozzle and a base material for collision therewith ofthe ejected particles; and pressure reducing means for reducing theinternal pressure of said sealed chamber.
 8. A thermal spraying systemcomprising: a thermal spraying nozzle device wherein carrier gas isintroduced from an inlet side of a nozzle to form a supersonic gas flowand a thermal spraying material is atomized and ejected by said gasflow, said thermal spraying nozzle device comprising, a storage sectionstoring molten metal as said thermal spraying material connected to anend on the inlet side of said nozzle through a connecting pipe, and,said nozzle having a throat portion and a divergent region in adownstream of said throat portion toward an outlet side to form thesupersonic gas flow, wherein said thermal spraying nozzle device isconfigured such that metal particles atomized by the supersonic gas floware cooled to a solidified or semi-solidified state in said divergentregion and then ejecting in a predetermined direction from the outletside of said nozzle: a molten metal supply unit connected to saidstorage section through a connecting pipe to supply molten metal underpressure continuously to the molten metal in the storage section; and abase material supply unit for continuous supply of said base material.